The Thermal Expansion Characteristics

of Stainless Steel Wel d Metal

Thermal expansion data are established to help in the proper selection of austenitic stainless steel filler metals to be used for welding dissimilar metal joints BY J. VV. ELMER, D. L. OLSON AND D. K. MATLOCK ABSTRACT. Thermal expansion coeffi-cients for 28 stainless steel welds of varying composition have been mea-sured. A graphic method of predicting the coefficient of thermal expansion (CTE) for stainless steel welds centered about the DeLong diagram has been prepared from these data. An overall description of the behavior of the CTE as a function of composition was accom-plished by supplementing the data with published thermal expansion data of all Fe-Ni-Cr type alloys. Lines of constant expansion were then mapped on the Fe-Ni-Cr ternary diagram and were sub-sequently transposed to the Schaeffler diagram. Using these diagrams, the CTE for a wide range of ferritic and austenitic alloys can be predicted. Residual delta ferrite in a stainless steel weld was shown to reduce the CTE of the duplex austenite-ferrite microstruc-ture from that of a ferrite-free micro-structure. Results predict the CTE of residual delta ferrite to be 15.0 ^ m/ m/ C over a 20 to 400C (68 to 752F) temper-ature range; in addition, wi th the Thomas composite theory of thermal expansion, the CTE of any stainless steel wel d con-taining delta ferrite can be calculated. A decrease in specific volume was shown to accompany the transformation of metastable delta ferrite to austenite and sigma phase. This contractive dilation results in a strain of 4.5 X 10~5 per per-cent ferrite that transforms. It should be emphasized that this strain will, in general, produce stresses that are tensile in nature across a weld joint. Furthermore, the total strain due to the transformation is determined by the initial ferrite content and the time that the joint is held at an elevated temperature. Introduction Thermal expansion is a fundamental material property which relates dimen-sional changes of a material wi th changes in temperature. A convenient measure of thermal expansion is the mean linear coefficient of thermal expansion (CTE) and is defined as: A L o ( T2- T, ) ., =/ A LX 1 \V Al (I) where Li and L2 are the specimen lengths at temperatures T, and T2 respectively, L0 is the initial specimen length and cem is the mean CTE. Fundamentally, the CTE mea-sures the amount of strain, AL/L0, which accompanies a material wi th a change in temperature, AT. When dissimilar materials are joined and restrained such as in a welding pro-cess, changes in temperature will allow stresses as high as the yield stress t o develop in the joint. For example, Dal-cher et al. (Ref. 1) predicted hoop stress-es up to 34 ksi (234 MPa) at the root of a typical dissimilar metal pipe wel d due to the CTE mismatch alone. The stresses decay as a function of position away from the interface so that far from the interface the CTE mismatch stresses are zero. The stress distribution is determined by the wel d joint design and weldment geometry, while the stress level depends on both the difference in CTE between the weld and base metals and the tem-perature change from the stress free temperature. The stress free temperature is the ref-erence temperature at which no CTE stresses exist. Typically, in the as-welded condition, this temperature will be close to the solidus of the weld metal. Howev-Based on paper presented at the 62nd A WS Annual Convention held in Cleveland, Ohio, during April 5-10, 1981. J. W. ELMER, D. L. OLSON and D. K. MAT-LOCK are with the Colorado School of Mines, Golden, Colorado. er, if a postweld heat treatment (PWHT) operation has been used, stress relief may be sufficient to reduce the stresses such that the PWHT temperature will be stress free. If welds of dissimilar metals are used at high operating temperatures, the thermally induced stresses can decay by creep mechanisms to low levels with time. The weldment at the operating temperature will then be essentially stress free. However, stresses will develop when the wel d is cooled to room tem-perature, a particular concern if thermal cycling is frequent. Selection of the proper filler metal to weld dissimilar metals that minimizes CTE mismatch stresses requires an under-standing of the thermal expansion prop-erties of the wel d metal. A compilation of published CTE data by Bennett (Ref. 2) for stainless steel welds and base metals was an early attempt to characterize these properties; however, little other informa-tion is available on the CTE behavior of weld metal. Stainless steel filler metals are frequent-ly used to weld austenitic and ferritic steels. Typically, the austenitic stainless steel has a CTE that is 30% to 40% greater than that of a ferritic steel. This CTE mismatch is frequently responsible for problems that occur in these type of joints. One such problem exists with the ferritic Cr-Mo steel to austenitic stainless steel weld transition joint commonly used in power plants (Ref. 3, 4). Numerous service failures resulting in costly down-time have occurred (Ref. 5) in these wel d joints, because of fusion line cracking at the ferritic weld interface when Type 310 Cb and other types of stainless steel filler metals were used. Thermal expansion mismatch has been sited (Ref. 3-6) as one of the major problems wi th this particular wel d. Some degree of success has been obtained in preventing or at least postponing the cracking problem by using ERNiCr-3 nick-WELDINC RESEARCH SUPPLEMENT | 293-s el-base filler metal which has a lower CTE than Type 310 Cb stainless steel filler, and has a CTE between that of the t wo base metals (Ref. 5). Another problem that exists is cracking of weld cladded material (Ref. 7). Austenitic stainless steels are clad on ferritic steels to provide corrosion resistance. Typically these welds cover large surface areas relative to other wel d-ing processes, and an understanding of how the CTE varies wi th dilution woul d be helpful in selecting proper filler met-als. Before discussing the specific aspects of this study, it is necessary to point out that the CTE of a material is dependent upon the prior thermomechanical pro-cessing history and the temperature range over which it is measured. For example, a ferrite-free wrought, stainless steel will have a different CTE than its welded counterpart containing residual delta ferrite, or a 36% nickel-iron alloy (Invar) will have different CTE's de-pending on the mechanical processing history and the temperature range of interest. It is, therefore, necessary to compare the CTE data for materials that have had the same metallurgical process-ing history welds, for example and whose CTE were measured over identical conditions. Otherwise, the correlations can only be considered approximate. In general, however, these correlations allow engineering trends to be estab-lished by means such as the plotting of isoexpansion lines on diagrams as a func-tion of composition, using data acquired from many sources. Experimental Procedures Welding and Specimen Preparation A three step weld alloying procedure, similar to that of Lippold and Savage (Ref. 8), was used to produce welds of desired compositions. This procedure is illus-trated in Fig. 1. First, a GMAW process was used to lay a bead on Vi in. thick (12.5 mm) Type 316 stainless steel plate using Y\b in. (1.6 mm) diameter Type 308, 309, 310, 312, 330 or 410 stainless steel filler metal wire. Welding grade argon-2% oxygen was used as the shielding gas, and the travel speed was varied from 11 ipm (4.65 mm/s) t o 28 ipm (11.9 mm/s) to produce several different welds from each filler metal. Next, the weld rein-forcement was removed by machining, leaving the base metal intact. This process left a strip of the GMAW deposited filler metal down the center of the base plate. In the final step, the alloying filler metal and the base metal were melted together using a CTAW pass. Welding grade argon was used as the shielding gas wi th a 2% thoriated tungsten electrode ground to a 90 deg included angle. The travel speed of this pass was varied from 2.5 ipm (1.07 mm/ s) to 3.1 ipm (1.32 mm/s) to provide different amounts of dilution of the CMA deposit wi th the Type 316 stainless steel base metal. The different GMAW deposition rates and GTAW dilution rates were used to produce a total of 28 welds of different compositions from the six filler metals. For each filler metal, a series of welds was produced which varied in composition from the base metal to the filler metal. The welding parameters and resulting chemical compositions of the test welds are presented in Table 1. The chemical compositions were measured using emis-sion spectroscopy and interstitial analy-sis. In Table 1, the three numbers of the specimen designation refer to the stain-less steel filler metal type, and the letters refer to high, medium, l ow or zero depo-sition rate (GMAW) and dilution rate (GTAW). For example, 308HL refers to 308 filler metal, high deposition (GMAW), low dilution (GTAW). Note that a high heat input was necessary to produce a weld bead of sufficient size, such that a specimen for dilatometric analysis could be removed. The all-weld-metal dilato-metric specimen machined from the cen-ter line of the weld bead as shown in Fig. IC measures 1.5 in. (38 mm) long and 0.15 in. (3.8 mm) in diameter. Testing Procedure Thermal expansion measurements were made using a horizontally posi-tioned vitreous silica push rod dilatome-ter. Dilation curves were recorded as specimen length change versus tempera-ture using an X-Y recorder. NBS standard reference materials SRM 736 (copper) and SRM 737 (tungsten) were used to calibrate the dilatometer. The dilatometer calibration curve was established using the tungsten standard. The copper stan-dard was then tested and from its cor-rected dilation curve the accuracy of the dilatometer was determined to be 1. 8% error in the mean CTE as deter-mined from 11 data points taken at 25C (45 F) intervals in the range from 275 to 525C (527 to 977F). All testing was conducted in an argon atmosphere to prevent oxidation of the specimen. The heating rate was con-trolled at 5.5C/mi n (10F/min) and cooling curves were recorded during a furnace cool at cooling rates less than 5C/ mi n (9F/min). The specimen tem-perature was recorded using a single type K thermocouple that was kept in contact with the center of the specimen during the test. The length change, accurate to 1 X 10- 4 in. (2.5 X 10~3 mm), was recorded using an LVDT mounted in a micrometer stage which was separated from the furnace wi th a water cooled barrier. The experimental thermal expansion data were corrected to obtain the true dilation curve using the established cali-bration curve. A microprocessor assem-bly equipped wi th a digitizer was used to analyze and fit the true dilation curve to a fourth order polynomial of the form: L = L0 + aT + bT2 4- cT3 4- dT4 (2) where specimen length, L, at tempera-ture, T, can be calculated from the origi-nal length, L0, and polynomial coefficients a, b, c, d. Both instantaneous and mean CTE data were calculated from this equa-tion. The polynomial estimation typically fit the actual data with a coefficient of GMA Bead 308,309,310,312,330,410 Filler 316 Stainless Steel 0.5 Specimen 0.15x (inch) GTA Pas s ipm Fig. 1 Three step weld alloying procedure used to produce stainless steel welds 294-s I SEPTEMBER 1982 Table 1Summary of Welding Procedures and Resulting Compos Specimen 308H* 308HL 308HH 309H* 309HL 309HM 309HH 309MM 309LH 31 OH* 310MM 310MH 310LH 330H* 330HL 330HM 330HH 330MM 410H* 410HL 410HM 410HH 410MH 410LH 312H* 312HH 312MM 312MH (a) I current in I 300 300 300 300 300 300 300 300 290 290 290 290 290 290 290 290 290 300 280 300 300 300 300 300 320 320 320 320 ampere* GMA<a> V 28-29 28-29 28-29 29-30 29-30 29-30 29-30 29-30 29-30 29-30 28-29 28-29 28-29 29-30 29-30 29-30 29-30 29-30 29-30 29-30 29-30 29-30 29-30 29-30 29-30 29-30 28-29 29-30 ; V voltage s 11 11 11 11 11 11 11 21 38 11 21 21 38 11 11 11 11 21 11 11 11 11 21 38 11 11 21 21 in volts i N/ A 310 310 N/ A 310 310 310 310 310 N/A 300 310 310 N/ A 305 310 310 310 N/ A 310 310 310 310 310 N/ A 310 310 310 S wel d GTA(a

V N/ A 16-17 16-17 N/ A 16-17 16-17 16-17 16-17 16-17 N/ A 17-18 17-18 17-18 N/ A 17-18 17-18 17-18 18-19 N/ A 18-19 18-19 18-19 18-19 18-19 N/ A 18-19 18-19 18-19 ng speed in s N/A 3.1 1.7 N/ A 3.1 2.7 2.1 2.7 2.1 N/ A 2.7 2.1 2.1 N/A 3.1 2.7 1.7 2.7 N/ A 3.1 2.7 1.7 1.7 1.7 N/ A 1.7 2.7 1.7 ipm. Cr 19.5 18.3 17.6 22.6 21.0 21.5 20.0 18.7 18.5 22.5 19.5 19.0 17.5 17.3 16.3 16.8 16.0 16.0 15.8 17.3 16.0 17.0 17.2 17.9 22.0 22.0 19.5 20.0 tion of 28 Stainless Steel Welds Si .38 .70 .78 .39 .44 .41 .36 .40 .42 .40 .40 .35 .39 .41 .33 .35 .35 .31 .37 .35 .32 .32 .32 .33 .50 .47 .45 .39 Mo 1.27 1.68 2.00 1.27 1.44 1.44 1.80 1.85 2.00 1.1 2.1 2.0 2.1 1.17 1.50 1.64 1.93 1.50 1.20 1.66 1.58 1.62 1.80 2.05 1.32 1.63 1.75 1.65 Nb Ti Zr <.01 <.01 <.01 <.01 <.01 <.01 <. 0I <.01 <. 0l <. 0l <.01 <.01 <.01 <. 0l <.01 < 0 1 <. 0I <. 0I <.01 <. 0I <. 0I <. 0I <.01 < 0 1 <. 0I <. 0I <.01 <.01 N .063 .034 .033 .059 .077 .100 .160 .100 .120 .075 .075 .078 .068 .042 .052 .079 .014 .079 .063 .100 .120 .220 .071 .075 .070 .079 .100 .080 Composition C .049 .052 .058 .071 .057 .056 .073 .060 .066 .082 .063 .065 .041 .013 .082 .010 .048 .075 .078 .072 .084 .072 .076 .072 .066 .058 .050 .076 Mn 1.34 1.49 1.53 1.53 1.72 1.80 1.33 1.37 1.25 1.35 1.35 1.30 1.35 1.54 1.47 1.43 1.30 1.22 .74 .90 .83 .87 .96 1.15 1.60 1.43 1.40 1.43 Ni 10.4 10.3 10.3 12.1 12.0 12.5 11.3 11.2 11.2 15.0 12.8 12.5 11.5 23.5 18.0 18.0 12.9 14.5 5.3 8.3 7.0 7.3 8.5 9.9 10.4 10.5 10.1 10.7 Cu .31 .09 .08 .27 .30 .31 .32 .39 .35 .23 .33 .32 .13 .09 .12 .11 .13 .11 .20 .28 .24 .25 .28 .30 .09 .11 .11 .24 Co .17 .21 .20 .19 .19 .20 .18 .18 .19 .12 .17 .18 .23 .13 .16 .15 .19 .17 10 .14 .12 .13 .15 .16 .18 .20 .20 .15 S .017 .005 .006 .024 .025 .025 .026 .027 .028 .016 .024 .025 .019 .019 .019 .018 .019 .019 .016 .022 .021 .021 .024 .025 .016 .018 .017 .024 O .054 .015 .013 .040 .020 .022 .029 .019 .023 .042 .052 .091 .045 .028 .039 .036 .033 .039 .030 .048 .042 .048 .049 .050 .041 .054 .044 .049 UJ 0 -O i UJ > LLI Q O OL < LU IA a. o - J > I o re < ui tn ui on I -z ul s D. o determination of 0.999999. The average error resulting from the digitizing and curve fitting is 0.50% as determined from the tungsten standard specimen. A magnetic inductive type ferrite meter was used to measure the ferrite content of the welds. Calibration of the meter was performed by comparing the magnetic ferrite measurements with point count measurements from t wo stainless steel wel d specimens, one con-taining 34.2% delta ferrite and the other containing 10.4% delta ferrite, and adjust-ing the amplifier gain on the meter to read identical with the point count mea-surements. During the investigation, it was deter-mined that the ferrite meter required correction when used to measure the ferrite content of the dilatometry speci-mens. Due to the specimens sharp radius of curvature, the volume-percent ferrite read by the meter was lower than the actual ferrite content. Therefore, the meter was calibrated by measuring the ferrite content of flat specimens and 0.15 in. (3.8 mm) diameter specimens taken from the same wel d. The resulting cali-bration curve, shown in Fig. 2, is pre-sented as a plot of ferrite content mea-sured on the curved surface vs. ferrite content measured on the flat specimen. All ferrite measurements reported in this paper represent an average of at least 10 data points taken on the 0.15 in. (3.8 mm) diameter dilatometry specimen and then converted to the actual ferrite content using the calibration curve of Fig. 2. Results and Discussion Thermal Expansion Coefficient and Chemical Composition A relationship between CTE and chem-ical composition can be used to predict 30 the thermal expansion properties of dis-similar metal welds by knowing the base metal compositions, filler metal composi-tion and the welding dilution. This rela-tionship was determined by experimen-tally measuring the CTE of various stain-less steel alloys and mapping this data as a function of composition. The DeLong a UJ oc 3 < UJ 2 UJ .2 C 20 0) E o a> a II) ^ *" 1 0 0C * a. "> UJ -^ o it? _ -1 1 1 FERRITE METER CALI BRATI ON CURVE , N V A8 f / / / / / / / / 1 1 A I / / A / yA / / / / / / / yA A X / / / / / I I I10 20 30 % FERRITE MEASURED, fl at speci men 40 > ul o I o re < ui Crt UJ re Q. O X < LU (rt Ul re CL. o > ui Q o re < ui v> ui re Fig. 2 Ferrite meter calibration curve to correct ferrite measurements made on the curved surface of the dilatometry specimens WELDING RESEARCH SUPPLEMENT | 295-s Table 2CTE, Creq, Specimen 308H$ 308HL 308HH 309H* 309HL 309HM 309HH 309MM 309LH 31 OH* 310MM 310MH 310LH 330H* 330HL 330HM 330HH 330MM 41 OH* 410HL 410HM 410HH 410MH 410LH 312H* 312HH 312MM 312MH Nieq and Ferrite Content for 28 Stainless Steel Welds Mean Coefficient Thermal Expansion, / i m/ m/ C, 20 18.0 18.5 18.4 17.3 18.2 18.3 18.5 18.9 18.6 17.7 18.3 18.1 18.7 18.9 18.4 18.4 18.6 I8.4 13.9 18.9 19.0 19.3 18.8 18.7 17.0 17.5 18.4 18.7 400 C Creq 21.3 21.0 20.8 24.5 23.1 23.6 22.3 21.2 21.1 24.2 22.2 21.5 20.1 19.1 18.3 19.0 18.5 18.0 17.6 19.5 18.1 19.1 19.5 20.5 24.1 24.3 22.2 21.9 Nieq 14.8 13.9 14.0 17.1 17.3 18.5 19.4 17.1 17.8 20.7 16.0 17.8 15.7 29.6 22.9 24.3 20.3 20.0 10.1 14.1 13.7 16.8 13.8 15.2 15.6 15.7 15.6 16.5 Ferrite, % 10.4 7.2 5.6 12.1 8.1 5.4 0.9 0.5 0.2 0 <0.2 <0.2 0.9 0 0 0 0 0 > 5 0w 4.0<a) 14(a) 1.6<a) 5.4<d> 3. 6( a ) 34.2 16.9 4.5 9.0 (a) Martensite is present. (Ref. 9), the Schaeffler (Ref. 10) and the Fe-Ni-Cr ternary diagrams were used to map the data as a function of austenite stabilizing elements (i.e., nickel equiva-lent) and ferrite stabilizing elements (i.e. chromium equivalent). The DeLong method of calculating the Ni equivalent and Cr equivalent: the amount that the CTE changes depends on the Creq to Nieq ratio. If the Creq is increased from this region, the composition enters the t wo phase A + F field. The residual delta ferrite in this region has a BCC crystalline structure and has a lower CTE than austenite. The influence that delta ferrite has on the CTE is discussed in detail in the next section and basically has the effect of reducing the CTE in a predictable manner. The weld wi th the highest ferrite con-tent, 312 H$ wi th 34.2% ferrite, has the lowest CTE of this region at 17.0 g.m/ m/ C; this represents a 8. 1% reduction in CTE from the austenitic base metal. The same effect in reducing the CTE woul d be expected if the composition were changed such that the A 4- F + M region was entered since martensite also has a lower CTE than austenite. One wel d containing a significant amount of mar-tensite, 410 H* , has a CTE of 13.9 gm/ m/ C thereby confirming this trend. This wel d, shown in Fig. 4, contains residual delta ferrite outlining the solidification structure in a martensitic matrix and has a hardness of Rc 37. Each series of filler metal compositions is indicated by a different symbol in Fig. 3 so that the effect of wel d dilution on the resulting composition and CTE can be followed. In comparing the CTE data for the welds from the 330 filler metal series and the 410 filler metal series, it is evident that changes in composition that result in formation of phases wi th thermal expan-sion properties different from the matrix being present have more effect on changing the CTE than do changes in composition that contribute only solid solution effects. For example, the 330 filler metal is high in nickel content and all the welds are fully austenitic with nickel equivalents in the range of 20 to 24. These welds have similar CTE of about 18.5 j t m/ m/ C. Since there are no other phases Nieq = Ni 4- 0.5 Mn 4- 30 (C 4- N) (3) (4) Creq = Cr 4- Mo 4- 1.5 Si was used. These equivalents, as well as the CTE and ferrite contents for the welds, are listed in Table 2. DeLong Diagram The majority of the welds have Cr and Ni equivalents that woul d place them in the compositional limits of the DeLong Constitution Diagram (Ref. 9). Figure 3 shows this diagram wi th the mean CTE plotted as a function of Cr and Ni equiv-alents. There are three different regions of interest on this diagram: the single phase austenite field (A), the t wo phase austenite and ferrite field (A 4- F), and the three phase austenite, ferrite and mar-tensite field (A 4- F 4- M). The CTE is influ-enced by the presence of these different phases. The highest thermal expansion coefficients were measured in the aus-tenite phase field at a composition of 14 Nieq and 18 Creq, where the CTE is around 19 ^ m/ m/ C. The CTE decreases as the composition is changed in any manner wi th respect to this point, and 18 20 22 24 CHROMIUM EQUIVALENT Fig. 3Mean coefficient thermal expansion (um/m/'C 20-400"C, i.e., 68-752F)plotted on the DeLong diagram as a function of chromium and nickel equivalents. Isoexpansion lines are included. Filler metals: 308, O 309, 310, A 312, J 330, V 410 296-s | SEPTEMBER 1982 Fig. 4 Optical micrograph sho wing martensit-ic stainless steel weld. Type 410 filler metal; Type 316 base metal; GMA weld. Fed} in HNOj etchant. X625 present to influence the CTE of fully austenitic welds, there should be a smooth transition in the CTE from 18.5 / * m/ m/ C to 15.5 / um/ m/ C as the nick-el content is increased to that of 100% nickel. Therefore, a large increase in nick-el content is necessary to reduce the CTE of a fully austenitic alloy by a significant amount. In contrast, the 410 filler metal series is different in the sense that a small compo-sition variation changes the CTE from 19 /t t m/m/C to values as low as 11 gm/ m/ C because of the presence of mar-tensite and ferrite in the microstructure. Thus the 410 GMA weld containing fer-rite and martensite has a CTE of 13.9 / t i m/ m/ C while the remaining welds in the 410 series, which were heavily dilut-ed with the Type 316 base metal, have CTE values around 19 f t m/ m/ C Fe-Ni-Cr Ternary and Schaeffler Diagram The narrow compositional range of the DeLong diagram limits its use to stainless steels. Therefore, to obtain an overall understanding of the variation in CTE wi th composition, CTE data was col-lected from the literature (Ref. 11-20) pertaining to all Fe-Ni-Cr containing alloys. Similar to the procedure used above wi th the DeLong diagram, the data were plotted on the ternary diagram and isoexpansion lines were drawn using these data as a guideline. CTE as a func-tion of composition for the Fe-Ni binary system (Ref. 17, 18) and for the Fe-Cr binary system (Ref. 19, 20) was particular-ly helpful in establishing the contours. Regions where the CTE data exists are shaded. Extrapolation of the contours outside these regions was accomplished by understanding the phases which are present (Ref. 12) and the CTE end points established from work on the binary systems (Ref. 17-20). The ternary diagram is shown in Fig. 5. The region of highest thermal expansion is again seen in the conventional stainless steel alloys centered about 14 Nieq and 18 Creq. This region represents a peak of CTE at approximately 19 nm/ m/ C. Reduction of the Nieq will result in delta ferrite being present and will decrease the CTE, while a reduction of both the Nieq and the Creq will result in delta ferrite and martensite being present and will also decrease the CTE. The lowest ther-mal expansion coefficients are observed in the Fe-36Ni, Invar-type, alloys; howev-er, these alloys are not frequently used for welding. Large variations in the CTE are appar-ent over the ternary diagram. The most significant changes that occur for ferritic-austenitic stainless steel welds are those that occur due to the presence of delta ferrite and martensite in the microstruc-ture. Therefore, the isoexpansion lines drawn on the ternary diagram were transposed to the conventional Schaef-fler Diagram (Ref. 10) so that both the constitution of the wel d and its CTE can be predicted from one diagram. This diagram is shown in Fig. 6. The DeLong diagram is outlined to illustrate where the initial data on welds is located, while the shaded areas again represent those regions where CTE data was compiled from the literature. This is a clearer repre-sentation of how ferrite and martensite influence the CTE. The martensite causes a trough of low expansion in the diagram while ferrite results in a more gradual decrease in CTE. The Influence of Delta Ferrite on the CTE Stainless steel welds will contain residu-al delta ferrite in the as-welded micro-structure if the composition of the weld has a sufficiently high Creq/ Nieq ratio. The residual delta ferrite has a BCC crystalline structure, and can be present in various morphologies (Ref. 21) within the FCC austenitic matrix. On solidification, there is a partitioning of ferrite stabilizing ele-ments to the delta ferrite and austenite stabilizing elements to the austenite matrix (Ref. 21). It has already been shown that solid solution effects on the CTE are small and that the important factor in determining the CTE of the material is the CTE of the phases that are present. To determine the CTE of residual delta ferrite and thus its contribution to the CTE of the duplex structure, a series of dilato-metric tests were conducted on a single specimen, 312H<J>, initially containing 34.2% delta ferrite. The residual delta ferrite is metastable and will transform to austenite and sigma phase at elevated temperatures. This phase transformation Cr Mean Coefficient of Thermal Expansion 20-400 C (/im/m)/C 4 6 8 Fig. 5 Mean coefficient thermal expansion (fim/m/C, 20-400C, i.e., 68-752 F) plotted as isoexpansion contours on the Fe-Nieq-Creq ternary diagram. Shaded areas represent regions where data were acquired H Z UJ _ l < > o UJ - J UJ tt o Mean Coef f i ci ent of Thermal Expansi on | j i m/m)/C, 20-400C 810 S^2l 4; i 6^17 Ferrite j I 8 16 24 32 CHROMI UM EQUI VALENT Fig. 6Mean coefficient thermal expansion (nm/m/C, 20-400C, i.e., 68-752F) plotted as isoexpansion contours on the Schaeffler constitution diagram. Shaded areas represent regions where data were acquired WELDING RESEARCH SUPPLEMENT | 297-s Table 3Ferrite Content, CTE, and Length Measurements After Each of Six Dilatometric Tests Conducted on a High Ferrite Weld Test no. vs wel ded 1 2 3 4 5 6 Maximum temperature of dilatometric test, C -550 650 700 750 800 800 Ferrite content after test vol % 34.2 32.7 26.4 18.0 7.0 3.1 2.0 Mean coefficient of thermal expansion, (Ai m/m)/C 2 0 ^ 17.2 17.4 17.5 17.8 18.1 18.2 -400C Specimen As determined from dilatometric measurements, in. 1.5068 1.5068 1.5058 1.5050 1.5048 1.5047 (mm) -(38.273) (38.273) (38.247) (38.227) (38.222) (38.219) length Micrometer measurements, in. 1.5070 1.5068 1.5062 1.5058 1.5051 1.5049 1.5048 (cm) (38.278) (38.273) (38.257) (38.247) (38.230) (38.224) (38.222) has been studied (Ref. 22-24), and the kinetics of this interface controlled diffu-sion transformation (Ref. 22) allow it to occur under the thermal cycling conditions of a dilatometric test. There-fore, the specimen was run through sev-eral dilatometric cycles that served t wo purposes: 1. To heat the specimen and trans-form only part of the ferrite. 2. To measure the CTE of the material, as a function of ferrite content. A separate metallographic specimen was placed in the dilatometer and sub-jected t o the same thermal cycling as the dilatometric specimen. A section of the metallographic specimen was removed after each thermal cycle to evaluate the microstructure at each ferrite level. Table 3 summarizes the testing proce-dures, resulting ferrite contents and CTE of this specimen. The initial dilatometric tests were run at lower maximum tem-peratures than the latter tests to limit and control the amount of ferrite that trans-formed. Figure 7 illustrates the CTE as a function of delta ferrite content. As expected, the CTE decreases wi th increasing ferrite content. The as-welded material containing 34.2% delta ferrite has a CTE of 17.2 f t m/ m/ C and after five thermal cycles the wel d contains 3.1% ferrite and has a CTE of 18.2 / i m/ m/ C. Since the residual delta ferrite is distrib-uted as a second phase throughout the austenite matrix, the CTE of the duplex structure may follow a composite theory behavior with delta ferrite content. A review of thermal expansion composite theory relationships by Nielsen (Ref. 25) compares three thermal expansion com-posite relationships for particulate filled systems. These relationships are the Kern-er, Thomas and Turner theories for ther-mal expansion. The Thomas equation was selected t o represent the expected behavior be-tween the CTE and delta ferrite content because, as Nielsen (Ref. 25) summarizes, the Thomas and the Kerner equations are more accurate in general than the Turner equation and because the Kerner equa-tion requires knowledge of the elastic properties which were not measured in this study. The Thomas theory states: In ctc = <t>F In <*F 4- <*>m In (5) where a is the CTE, $ is the volume fraction, and the subscripts C, M and F refer to composite, matrix and fiber respectively. A regression analysis was performed, fitting the data to the mathematical form of the Thomas equation. Extrapolation of this regression line to 0% ferrite predicts the CTE of a fully austenitic material of this composition to be 18.4 / um/ m/ C and enables the CTE of ferrite to be calculated from the Thomas equation. The CTE of the residual delta ferrite as determined by this analysis is 15.0 gm/ m/ C. Wi th the CTE of the residual delta ferrite, the Thomas theory can be used to calculate the CTE of a stainless steel wel d containing delta ferrite. The CTE of the fully austenitic material can be approxi-mated by the CTE of the stainless steel base metal or filler metal of the same approximate composition in the ferrite free condition, the ferrite content can be measured, and the CTE can be calcu-lated. For example, if an autogeneous weld is made on Type 304L stainless steel base metal resulting in a residual ferrite content of 8%, the CTE of the wel d is calculated from equation (5), with the CTE of the austenitic phase taken as the CTE for wrought Type 304 stainless steel material (18.2 f t m/ m/0C, Ref. 12), to be 17.9 / xm/ m/ C. Figure 8 compares the initial and final microstructures of the thermally cycled Type 312 stainless steel wel d. Figure 8A shows the as-welded microstructure; the dark etching delta ferrite is present in a continuous vermicular morphology fre-quently observed in stainless steel welds. After three thermal cycles (Fig. 8B), the ferrite (dark) has transformed to sigma phase (black) and austenite (light). Posi-tive confirmation of the three phases was o (A z < a X Ul -1 < o 5 " f 0 Uj o I * "-o h- CM z UJ O 0 \ iL fc u. v UJ fc O i o z < UJ z 18. 5 1b 1 7 1 7 0 5 0 DELTA FERRI TE, Vol ume % Fig. 7 CTE plotted as a function of residual delta ferrite content illustrating the reduction of CTE with increasing delta ferrite content 298-s | SEPTEMBER 1982 Fig. 8 Optical micrograph showing the microstructure of Type 312 stainless steel weld in: Aas-welded condition; B thermally cycled condition. KOH electrolytic etch. X625 accomplished using the magnetic etching technique of Gray, Sikka, and King (Ref. 26) and using Murakamai's etch (Ref. 26); however, a 10% KOH electrolytic etch (1 A/ cm2 for 5 to 7 s) produced the same results and was less time-consuming. Figure 9 shows the same microstruc-ture at a higher magnification. Figure 9A shows a region of the microstructure where the delta ferrite transformed to sigma phase and austenite in approxi-mately equal amounts; it is to be noted that there appears to be a crystallograph-ic orientation relationship between the newly formed sigma phase and austenite and the original delta ferrite. The transfor-mation of delta ferrite was not uniform throughout the microstructure. Figure 9B shows a region of the same specimen where much less of the delta ferrite transformed and the sigma phase that is present is not as abundant as that of Fig. 9A. The original ferrite-austenite boundaries are decorated with what seem to be precipitates and they indicate the size of the delta ferrite particle prior to thermal cycling. After six thermal cycles, a point count of sigma phase shows 14.9% to be present in the microstructure and ferrite measurements show 2.0% ferrite left untransformed. This means that of the initial 34.2% ferrite, only 17.3% of the ferrite in the microstructure transformed directly to austenite. The presence of sigma phase in the microstructure and its influence on the CTE has not been considered up to this point. However, if the regression line of Fig. 7 is extrapolated to 0% ferrite, wi th 14.9% sigma phase in the austenitic struc-ture, the material has a CTE of 18.4 / um/ m/ C. Comparing this value to that of about 18.5 ^ m/ m/ C for the fully austenitic structures shown in Fig. 3 indi-cates that there is little difference in the CTE, although one structure contains 14.9% sigma phase and the other is fully austenitic. This leads to the conclusion that sigma phase has a CTE that is nearly the same as that of austenite. To account for sigma phase, a separate term could be included in equation (3); however, the extra term is not necessary to determine first order effects since the effective CTE of sigma phase in austenite is close to that of the austenitic matrix. Decrease in Specific Volume as Delta Ferrite Transforms to Austenite Irreversible dilatometric behavior was observed between heating and cooling cycles of the welded specimens when: 1. The maximum temperature of the test went above about 600C (1112F). 2. The weld contained delta ferrite. In all cases where irreversible behavior was observed, the cooling curve woul d lie below the heating curve. This type of dilatometric behavior of stainless steel welds has also been observed by Bloch and Huszar (Ref. 24) and is associated wi th the transformation of delta ferrite. The difference between the heating and cooling curves represents a contraction of the specimen, independent of the thermal expansion (contraction) proper-ties of the material. This behavior is illus-trated in Fig. 10, which compares t wo dilatometric curves of the high ferrite wel d series described in the previous section. Figure 10A shows the typical dilation behavior of the wel d when 11.0% of the ferrite transforms. The heating and cool-ing curves do not superimpose; in fact, the cooling curve lies significantly below the heating curve. There is a negative displacement in the specimen length of 0.0008 in. (0.02 mm), measured at room temperature after the test; this means that the specimen has contracted. In contrast, Fig. 10B shows that the heating and cooling curves almost superimpose when only 1.1% of the ferrite transforms. Wi th subsequent thermal cycling, the material woul d stabilize and the dilato-metric curve woul d be reversible. The change in specimen length is due to a decrease in specific volume of the material as the delta ferrite transforms to the more dense austenite and less dense sigma phase. To estimate how much contraction should occur, the densities of Type 410 ferritic stainless steel (7.7 gm/ cm3 Ref. 13), Type 316 stainless steel (8.0 gm/ cm3- Ref . 13), and Fe-Cr sigma phase (7.6 gm/ cm3 Ref. 27) were used to approximate the densities of the phases present in this wel d. Wi th these density approximations, a welded material containing 34.2% ferrite and 65.8% austenite woul d have a densi-ty of 7.90 gm/ cm3, while the same mate-rial after thermal cycling containing 2% ferrite, 83.1% austenite and 14.9% sigma phase woul d have a density of 7.93 gm/ cm3. The difference in specific vol-ume between the material in the as-welded and thermally cycled conditions woul d be 0.378%. The relationship, AV = 3em, (Ref. 28) where AV is the change in specific vol -ume and tm is the linear strain, is good for small strains and can be used to calculate the strain in any direction to be 1.26 X 10~3 and for the 32.2% change in ferrite, there should be 3.9 X 10~5 strain for each percent ferrite that transforms. Note that this strain woul d be greater if less sigma phase had been formed as would be the case if a stainless steel base metal less susceptible to forming sigma phase than Type 316 stainless steel woul d have been used. This value of strain can now be compared to the measured val-ue. Specimen length as a function of ferrite content is summarized in Table 3 and plotted in Fig. 11. The specimen length at each ferrite level was determined by subtracting the displacement, measured from the dilatometric curve at room tem-perature, from the previously calculated length. The specimen length was also measured with a micrometer after each test and both measurements are included on this plot. <r Fig. 9 Optical micrograph of thermally cycled Type 312 stainless steel weld showing: A region where delta ferrite transformed to sig-ma phase and austenite in equal amounts; B region where very little sigma phase is present upon delta ferrite transformation. KOH electrolytic etch. XI500 (reduced 32% on reproduction) WELDING RESEARCH SUPPLEMENT | 299-s = 20 E H O z LU J 1 0 -UJ o z < Dilatometric Test #3 312 Filler = 20 E O z UJ UJ o z < IU 10 0 i i Dilatometric Test # 6 312 Filler JT i > /L= 1. 5049 in . 80= 3 . 1 % 8F= 2 . 0 % i i i 800 22 200 4 0 0 6 0 0 800 TEMPERATURE, C TEMPERATURE, C Fig. 10 - Dilatometric curves illustrating a dilation of the specimen as delta ferrite transforms to austenite and sigma phase: A - transformation of 11.0% delta ferrite; B transformation of 1.1% delta ferrite The specimen in the as-welded condi-tion was 1.5070 in. (38.278 mm) long, and after six thermal cycles the specimen had contracted to 1.5048 in. (38.222 mm). Thus, the specimen decreased 0.0022 in. (0.056 mm) in length with a 32.2% decrease in ferrite. The change in speci-men length wi th ferrite content can be represented by a linear relationship; this is expected since the volume strain is a direct function of the linear strain for isotropic behavior. The change in specimen length is the strain which accompanies the phase transformation of delta ferrite to austen-ite and sigma phase and from this data was calculated to be 4.5 X 10- 5 strain for each percent ferrite that transforms. The estimated value for strain, 3.9 X 10- 5 for each percent ferrite that transforms, is lower than the measured value but con-sidering the approximations made to obtain the estimated value there is a good correlation. The point to be made, however, is that there is a theoretical basis for the observed contraction. The implication of a decrease in specif-ic volume during the phase transforma-tion is that, if a stainless steel weld con-tains delta ferrite and is held at an ele-vated temperature such that transforma-tion of the ferrite occurs, stresses will develop in the wel d joint that will be tensile in nature. These tensile stresses will be over and above the conventional CTE mismatch stresses that are present in the dissimilar metal wel d. CTE mismatch stresses, stress assisted oxidation at the interface, and acceler-ated creep in a carbon depleted zone close to the wel d interface have been cited as problems contributing to the failures of austenitic-ferritic dissimilar met-al welds (Ref. 3-6). Stress state at the wel d interface is the source of these problems and thus the source of the subsequent cracking that occurs. Thermal expansion mismatch stresses, operating stresses, and static loading stresses contribute to the high stress state at the interface, and the stresses resulting from the phase transformation must also be considered for those welds that contain delta ferrite and are operated at high temperatures. Because the transformation of delta fer-rite is an activated process, the stresses that develop from the transformation will do so as a function of time at operating temperatures. Therefore, the stresses in the wel d joint may develop over a period of time. Conclusion In summary, a means t o predict the coefficient of thermal expansion of stain-less steel welds has been developed by plotting isoexpansion lines on the DeLong constitution diagram. Twenty-eight data points taken from stainless steel welds provided the means to plot the isoexpan-sion lines. In addition, the CTE data, supplemented by literature data, was plotted on the Fe-Ni-Cr ternary diagram and on the Schaeffler constitution dia-gram for wel d, wrought and cast materials. Isoexpansion lines were plotted on these diagrams to estimate the CTE as a function of composition; however, these lines can only be considered approximate since the CTE can vary with material processing history. The conclu-sions from this portion of the study are: Changes in composition that result 1. 508 1 . 5 0 4 312 Filler metal o- Dilatometer a Micrometer 10 20 30 DELTA FERRI TE, Vol ume % 40 Fig. 11- Specimen length vs. delta ferrite content showing the contraction of the specimen as delta ferrite transforms to austenite and sigma phase 300-s | SEPTEMBER 1982 in mart ensi t e or f erri t e bei ng present in t he mi cr ost r uct ur e have t he largest ef f ect on t he CTE of t he stainless steel wel d. Ferrite f r ee stainless steel al l oys have t he hi ghest CTE of Fe-Ni-Cr t er nar y al l oys. This maxi mum in CTE occur s at a c o mp o -si ti on of 14 Ni equi val ent , 18 Cr equi va-l ent, and has a val ue of about 19 / xm/ m/ C, 20- 400C ( 68- 752F) . The CTE is rel at i vel y i nsensi ti ve t o changes in composi t i on t hat occur in t he single phase austeni te f i el d whi ch cont r i b-ut e onl y sol i d sol ut i on ef f ect s t o t he CTE. Ferrite present in a stainless steel wel d wi l l r educe its CTE f r om t hat in t he ful l y austeni ti c state. The r educt i on in CTE cor r esponds t o somet hi ng less t han 10% f or t ypi cal stainless steel wel ds. The mean CTE of resi dual del t a f er r i t e in a dupl ex st ruct ure was cal cul at ed t o be 15. 0 yum/ m/ C, 20- 400C ( 68- 752F) , and con-t ri but es t o t he CTE of t he dupl ex struc-t ur e by t he Thomas t heor y of t her mal expansi on f or composi t e materi al s. Usi ng this equat i on and t he CTE of resi dual f er r i t e, t he CTE of a stainless steel wel d can be cal cul at ed knowi ng t he del t a f er-ri te cont ent . A di l ati on is associ ated wi t h t he trans-f or mat i on of met ast abl e del t a f erri t e t o austeni te in f erri t e cont ai ni ng stainless steels. This t r ansf or mat i on occurs at el e-vat ed t emper at ur es as a f unct i on of t i me and cor r esponds t o a decrease in t he speci fi c vol ume of the mat eri al . A con-t ract i ve strai n of 4.5 X 10~5 was mea-sured and s hown t o accompany each per per cent f er r i t e t hat t ransf orms t o aust en-ite and si gma phase. This strai n is di rect l y associ ated wi t h this phase t r ansf or mat i on and can have t he ef f ect of l oadi ng t he wel d j oi nt i n t ensi on. Acknowledgment The aut hors wi sh t o acknowl edge t he research suppor t of t he Uni t ed States Depar t ment of Energy and t he assistance of t he Republ i c Steel Cor por at i on. Tech-nical discussions wi t h G. D. Ries of t he Republ i c Steel Cor por at i on and T. A. Whi ppl e of t he Nat i onal Bureau of Stan-dards wer e hel pf ul and are grat ef ul l y appr eci at ed. References 1. Dalcher, A. W Yang, T. M., and Chu, C. L. 1977. High temperature thermal-elastic analysis of dissimilar metal transition joints. Journal of Engineering Materials and Technolo-gy 99 (1):65-69. 2. Bennett, A. P. 1969 (December). 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Transactions of the ASME, Journal of Materials Engineering and Technology 98(4): 348-356. 8. Lippold, ). C, and Savage, W. F. 1980. Solidification of austenitic stainless steel wel d-ments: part 2 the effect of alloy composition on ferrite morphology. Welding journal 59(2): 48-s to 58-s. 9. DeLong, W. T, Ostrom, C. A., and Szumachowski, E. R. 1956. Measurement and calculation of ferrite in stainless-steel wel d metal. Welding Journal 35(11): 521-s to 528-s. 10. Schaeffler, A. L. 1949. Constitution dia-gram for stainless steel wel d metal. Metals Progress 56: 680. 11. Touloukian, Y. S., Kirby, R. K., Taylor, R. E and Desai, P. D. 1975. Thermophysical Properties of Matter Volume 12 Thermal Expansion, New York-Washington; 1FI, Ple-num. 12. Peckner, D., and Bernstein, I. M. 1977. Handbook of stainless steels, McGraw Hill. 13. The International Nickel Company, Inc. 1968 (November), Properties of some metals and alloys. 3rd ed. 14. Climax Mol ybdenum Company. 1971. 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The alloys of iron and chromium, volume IIhigh chro-mium alloys, p. 93. New York and London: McGraw Hill, Inc. 21. David, S. A. 1981. Ferrite morphol ogy and variations in ferrite content in austenitic stainless steel welds. Welding Journal 60 (4): 63-s to 73-s. 22. Raghunathan, V. S., Seetharaman, V., Venkadesan, S., and Rodriguez, P. 1979 (No-vember). The influence of post wel d heat treatments on the structure, composition and the amount of ferrite in type 316 stainless steel welds. Metallurgical Transactions A, 10A: 1683-1689. 23. Wegrzyn, (., and Klimpel, A. 1981. The effect of alloying elements on sigma phase formation in 18-8 weld metals. Welding lour-nal 60(8): 146-s to 154-s. 24. Bloch, R., and Huszar, R. 1980. The disintegration of delta ferrite in welding mate-rial X 3 CrNi MoN 25 20. Metallography con-ference, Augsburg (GFR). 25. Nielsen, L. E. 1967. Mechanical proper-ties of particulate-filled systems, journal Com-posite Materials I: 100-119. 26. Gray, R. I., Sikka, V. K., and King, R. T. 1978. Detecting transformation of delta-ferrite to sigma-phase in stainless steels by advanced metallographic techniques. /. Metals 30(11): 18-25. 27. Duwez, P., and Baen, S. R. 1950 (June). X-ray study of the sigma phase in various alloy systems. ASTM STP No. 110: 48-60. 28. Dieter, G. E. 1976. Mechanical metallur-gy, 2nd ed., p. 44. Mcgraw-Hill. WELDING RESEARCH SUPPLEMENT 1301-s